Coal Energy Conversion with Aquifer-Based Carbon Sequestration:
An Approach to Electric Power Generation with
Zero Matter Release to the Atmosphere
Final Report
Principal Investigators
Reginald E. Mitchell, Professor, Mechanical Engineering; Christopher Edwards,
Associate Professor, Mechanical Engineering; Scott Fendorf, Professor, Department of
Environmental Earth System Sciences
Post-doctoral and Graduate Student Researchers
Ben Kocar, Adam Berger, Eli Goldstein, John (J.R.) Heberle, BumJick Kim, Andrew
Lee, Paul Mobley, and Rebecca Pass
Abstract
Decarbonization of electricity production is a vital component in meeting stringent
emissions targets aimed at curbing the effects of global climate change. Most projected
pathways toward meeting those targets include a large contribution from carbon capture
and storage. Many capture technologies impose a large energy penalty to separate and
compress carbon dioxide (CO2). Also, injected neat CO2 in a deep saline aquifer is
buoyant compared to the aquifer brine and requires an impermeable seal to prevent it
from escaping the aquifer. The proposed alternative technology processes coal in
supercritical water pumped from a saline aquifer. The entire effluent stream is
sequestered, capturing all carbon and non-mineral coal combustion products in the
process. This stream is denser than the aquifer brine and therefore offers a higher level of
storage security, and could utilize aquifers without suitable structural trapping. This
technology also increases energy security in the United States, allowing for the use of its
coal resources while avoiding atmospheric pollution.
The overall objective of this project was to provide the information needed to enable
development of this technology that takes water from a deep saline aquifer, desalinates
and uses it to process coal at supercritical water (SCW) conditions, and returns the water,
which will contain dissolved CO2 and other combustion products, back to the aquifer.
The sensible energy of the hot combustion products is transferred to the working fluid of
a heat engine for electricity generation. The only system effluents are supercritical water
containing dissolved coal conversion products, and solids, composed of ash and salts that
precipitate from the supercritical water.
A complete architecture employing supercritical water oxidation has been studied,
including a liquid-oxygen-pumped air separation unit and regenerator system that heats
and desalinates the incoming brine. A system-level, thermodynamic analysis indicates an
overall plant thermal efficiency of 36.7%, taking into account all separation and storage
energy penalties. In addition, an exergy analysis gives insights into the least efficient
parts of the proposed system. Results from the model and these analyses elucidate how
the proposed system may be operated as a zero-emission electricity source and the
technical challenges that must be addressed for deployment.
A software library for the thermodynamic properties of CO2 and H2O, gases that
behave non-ideally at the temperatures and pressures employed in the process units, was
developed and used to determine properties in the thermodynamic analysis. In order to
determine the thermodynamic properties of mixtures, a mixing model was also developed
that models the departure from an ideal Helmholtz solution using a modified
corresponding-states approach. The model was used to determine thermodynamic
properties when predicting the composition of the synthesis gas produced in the coal
reformer and the composition of the combustion products.
Geochemical models were used to assess the fates of potentially toxic trace elements
in the coal that may be dissolved in the water being returned to the aquifer. Results
suggest that relatively high concentrations of As, Cr, Cu, Hg, Pb, Zn, and Cd in the brine
being returned to the aquifer are likely to be mitigated by secondary
precipitation/adsorption reactions that occur within aquifers. The extent of these
reactions will depend on the mineralogical makeup of the aquifer. At the low
concentrations of dissolved trace elements in the brine being returned to the aquifer, no
adverse impacts on aquifer ecology are forecast.
Experimental facilities were built to permit the study of the coal reformation process
in SCW environments as well as to permit the study of combustion of the synthesis gas
produced. The coal reformation study was designed to determine coal conversion rates to
synthesis gas and char over selected ranges of temperature and pressure near the critical
point of water (Tc,H2O = 647 K, Pc,H2O = 221 bar). A 15-meter long reactor that provides
up to 15 minutes of residence time at specified temperature and total pressure was built
for this purpose. A 50 kW facility was also built to demonstrate stable combustion of the
synthesis gas in supercritical water with near-stoichiometric amounts of oxidizer. Stable
combustion has been demonstrated recently, and further combustion studies are
underway.
Wyodak coal, a low-ash, low-sulfur coal, was selected as the base-case coal for the
experimental study of the reformation process in SCW. This is a sub-bituminous coal
from Wyoming. It is typical of the coals from the Powder River Basin that are currently
being used for electric power generation in the United States.
Introduction
This project had the objective of providing the information needed to develop a coal-
based electric power generation process that involves coal conversion in supercritical
water with CO2 capture and storage in an inherently stable form. The only process
effluents are supercritical water (containing dissolved coal conversion products), and
solids, composed of fly ash and salts that precipitated from the water. Having no gaseous
emissions, such a system eliminates the need for costly gas cleanup equipment. In
addition, it provides a carbon capture and storage (CCS) option that may be more
acceptable to the public than other proposed CCS options.
Carbon capture and storage is a class of approaches to allow for the burning of fossil
fuels with minimal emissions of CO2 to the atmosphere. Models projecting technology
paths toward meeting emission targets typically include a large CO2 reduction from CCS
[1]. While enhanced oil recovery and enhanced coal bed methane recovery offer value-
added operation of CO2 storage, deep saline aquifers offer the largest storage potential of
CO2 [2]. The CO2 is less dense than the aquifer brine in an underground reservoir.
Therefore, storage in saline aquifers requires a suitable stratigraphic trap (such as a
caprock) above the reservoir to prevent the CO2 from escaping back to the surface. Over
time the CO2 becomes more secure in the reservoir by residual trapping in the aquifer
pores, dissolving into the brine, and finally forming carbonate minerals.
The purpose of this research is to analyze the feasibility of a plant that bypasses the
need for stratigraphic and residual trapping by injecting the CO2 into the aquifer already
dissolved into the brine. Once dissolved in the brine, it is denser than the surrounding
fluid and will sink to the bottom of the reservoir rather than escape back to the surface.
This type of CCS could allow for decreased monitoring of injectant plumes and less
potential remediation. This dissolved injection could be accomplished using a
conventional separation system on a pulverized coal (PC) power plant [3]. One could
pump the brine to the surface, dissolve the CO2 in it, and pump the brine—and the
captured CO2—back underground. This process would require multiple separation
processes that each carry an energy penalty, reducing overall efficiency. Another method
to achieve CO2 dissolution in aquifer brine, which is the subject of this research,
incorporates a supercritical water oxidation (SCWO) power plant that pumps aquifer
brine to the surface, reforms and combusts coal in supercritical water generated from the
brine, and returns the combustion products in solution to the aquifer. Constructing power
plants that have no gaseous emissions at all but instead, that sequester the entire effluent
stream would end the expensive cycle of continually identifying and managing the next-
most-harmful coal conversion product.
Deep saline aquifers have been recognized as suitable locations for the storage of
CO2. Identified sites across the globe have an estimated capacity to store a minimum of
125 years of CO2 generated in coal-fired power plants at the rate of coal consumption for
electric power generation in 2005 [2]. A case study on a representative aquifer with 100
mD permeability and 12% porosity determined the effluent of a 500 MWe power plant
with a 50 year injection period would cover an area of 63 square miles [3].
The advantages of this aquifer-based, coal-fired power plant relative to current and
other proposed power generation systems include (i) maximally efficient power
production while storing CO2 products in indefinitely stable forms, (ii) zero traditional air
pollutant emission and stack elimination, and (iii) size reduction of reactor vessel
(compared to pulverized-coal systems). Although the thrust of the project is directed to
clean coal utilization, the process being developed applies in general to all stationary
thermal processors (gasifiers and combustors) using all types of fuel (coal, natural gas,
oil, biomass, waste, etc.). Our investigation aims to lay the foundation for an efficient
coal energy option with no matter release to the atmosphere and in which all fluid
combustion products, particularly carbon dioxide, are pre-equilibrated in aquifer water
before injection into the subsurface.
There have been few external developments in supercritical water oxidation during
the course of the research project. The most mature experimental SCWO work has been
performed in Switzerland with the goal of use in aqueous organic waste destruction [4].
The SCWO flame would provide an internal heat source for waste destruction processes.
The most recent work from 2007 was interested in minimizing reactor fouling and
plugging from precipitation of salts from the simulated waste fluid [5]. Recently, a low
emission power plant utilizing SCWO combustion of fine coal particles was proposed
and a system model was analyzed with a net efficiency of 28% [6]. This plant
incorporates a SCWO combustor with coal used directly as the fuel—not reformed—and
separates ash and CO2 from the effluent stream after combustion, in contrast to this
research project. Such an architecture does not avoid the separation energy penalties that
this research aims to circumvent, leading to a lower efficiency than the architecture
proposed below.
The Proposed Process
In the proposed energy conversion scheme, a coal slurry, oxygen and water obtained
from a deep saline aquifer are fed to a gasification reactor (a reformer) that is maintained
at conditions above the critical temperature and pressure of water. Due to the solvation
properties of supercritical water, the small polar and nonpolar organic compounds
released during coal extraction and devolatilization are dissolved in the water and the
larger ones are hydrolyzed, yielding dissolved H2, CO, CO2, and low molecular weight
hydrocarbons, without tar, soot or polyaromatic hydrocarbon formation. In essence, a
synthesis gas dissolved in SCW is produced in the reformer. The syngas is sent through a
solids separator in which the solids are returned to the reactor (and ultimately separated),
and the syngas is directed to a combustor where it reacts with oxygen to produce
combustion products dissolved in supercritical water. The process stream that exits the
combustor is a single-phase, supercritical solution of hot combustion products. The
stream is passed through a heat exchanger, transferring its sensible energy to the working
fluid of a heat engine that produces electrical power in a combined-cycle scheme. The
cooled water stream, saturated with combustion products, is returned to the same or
nearby aquifer.
Since inorganic salts are insoluble in supercritical water, the aquifer water must be
desalinated before use. If not, the ability of the water to absorb coal conversion products
would be reduced and the salts would precipitate in the gasification reactor, mixing with
the ash. The salts would have to be separated from the ash if the ash were to be used, for
instance, as aggregate material.
Sulfur, nitrogen and many of the trace elements in coals are converted to insoluble
salts in SCW. The salts formed will precipitate from the fluid mixture in the reformer,
and will be removed from the system with the coal ash. In this coal conversion scheme,
all trace species introduced with the coal (such as mercury, arsenic, etc.), as well as all
coal conversion products, are sequestered in the aquifer along with the CO2. The only
matter not directed to the aquifer is the solid material consisting of the ash and
precipitated inorganic salts. Although the proposed system still has an energy cost of air
separation, the trade-off of carbon separation for air separation is advantageous since in
the process all fluid coal combustion products are sequestered, including sulfur and trace
metals, not just carbon dioxide.
Research Objectives and Tasks
The overall objective of this research project was to provide the information needed
to design and develop the key process units in the proposed aquifer-based coal-to-
electricity power plant with CO2 capture and sequestration. The project was divided into
four research areas - Area 1: Systems Analysis, Area 2: Supercritical Coal Reforming,
Area 3: Synthesis Fluid Oxidation and Heat Extraction, and Area 4: Aquifer Interactions.
In Area 1, thermodynamic analysis of the scheme provided fully qualified cycle
efficiency and process analyses. These analyses were used to determine the design
choices made in investigating component requirements in the proposed scheme.
Research efforts in Area 2 (Supercritical Coal Reforming) were aimed at determining
the supercritical water conditions that maximize the amount of chemical energy from the
coal in the synthesis fluid. Defining the optimum amount of oxygen required to drive the
gasification reactions and at the same time yield a high energy-content synthesis fluid as
a function of coal composition in the SCW environment was one of the goals of this task.
In concert with this was the goal of developing models that can predict accurately coal
conversion rates to synthesis fluid under SCW conditions. This required obtaining the
data needed to characterize coal extraction and pyrolysis rates and char gasification and
oxidation rates in supercritical water environments as functions of temperature, pressure
and properties of the coal and its char.
The research efforts associated with Area 3 (Synthesis Fluid Oxidation and Heat
Extraction) were focused on the design of the oxidation reactor and transfer of the energy
released to a heat engine in order to extract work. The stream exiting the oxidation
reactor, entering the heat exchanger of the heat engine needs to operate as close as
possible to material thermal limits to maximize heat engine efficiency. Thus, a primary
goal of the oxidation reactor design effort was the distancing of oxidation zones from
reactor walls. An additional requirement was control of reducing and oxidizing streams
to avoid liner corrosion. Under consideration was the design of a combustor in which
hydrodynamics and water injection were used to control reaction, mixing, and wall
interactions.
Research activities associated with Area 4 (Aquifer Interactions) were concerned with
characterizing the impact of dissolved constituents in the water being returned to the
aquifer on aquifer ecology. Of interests were the fates of contaminants prevalent within
coal, such as arsenic, mercury, and lead. Geochemical conditions in the deep subsurface
are likely to lead to the partitioning of elements such as As and Hg to the solid phase.
These elements are subject to migration should physical isolation be disturbed. Another
concern was the possible oxygenation of the aquifer, potentially destabilizing the sulfidic
minerals. A third concern was the potential to develop dramatic fluctuations in pH
resulting from variations in CO2 content, possibly destabilizing aquifer solids and
inducing dissolution or colloidal transport. Geochemical constraints are expected to
diminish the risk imposed by heavy metal discharge into the physically isolated deep
brines but in the research efforts, a combination of equilibrium based predictions and
spectroscopic/microscopic characterization of the energy system products were
performed to verify reaction end-points.
Materials degradation represents one of the most critical issues in the development of
the proposed process. The simultaneous presence of oxygen and ions in the supercritical
fluid forms an aggressively corrosive environment. Consequently, reactor designs must
minimize the contact between walls and SCW that contains oxygen. In our approach, a
perforated liner was considered for use inside the combustor that permits cooling flows of
pure water to protect the combustor surfaces from the hot oxygen-containing SCW.
In the following sections, significant results of each Area are discussed.
Area 1: Systems Analysis (Edwards)
System Model
The proposed system model has evolved from its initial concept following the
detailed analysis of the plant. Initially, a single regenerator/desalinator heat exchanger
was employed that preheated the aquifer brine and oxygen with the combustion products
after they exited the main heat exchanger of the heat engine. The two processes are now
split into separate systems, with an active desalinator operated as a two-pressure-level
system. This change avoids difficult practical implications brought on by handling salt
precipitation in a single regenerator/desalinator as the brine becomes supercritical [7].
This also avoids a pinch point that results from the difference between pseudocritical
points in the two streams.
The ASU is modeled as a liquid oxygen (LOX) plant with a high pressure pump.
The LOX pump allows for increasingly high pressures to be achieved with minimal
change in the work investment. The SCWO system is not analyzed in detail here because
the system can be modeled thermodynamically without regard to internal details. The
regenerator/desalinator is discussed in detail below, as well as the heat engine. The ASU
was analyzed completely, but is only discussed briefly in this report.
Aspen Plus 7.1 was used to analyze the system model. A mixed set of property
methods were used throughout the system including the NBS steam tables for water in
the Rankine cycle and after desalination (until entry into the SCWO system). The
electrolyte-NRTL method was used for the aquifer brine through desalination. The Peng-
Robinson property method with Boston-Mathias extrapolations for high temperature
gases was utilized for the remainder of the system.
Figure 1 shows the interactions of the main plant components. The process begins
with the aquifer brine pumped to the surface and entering the regenerator/desalinator.
About 10% of the brine is desalinated, since only that portion is required by the SCWO
system to moderate the combustion process. The remaining brine is needed to keep all of
the combustion products, including CO2, in solution at re-injection. The amount of brine
necessary for dissolution is calculated from published CO2 solubility relations [8]. Neat
oxygen also enters the regenerator/desalinator from an ASU and is preheated. Air
separation is necessary because the solubility of nitrogen in water is much less than that
of CO2, and its inclusion would prohibitively increase the amount of brine needed to
dissolve the combustion products. Some of the desalinated water goes to coal slurry
preparation. The slurry is sent back to the regenerator/desalinator to be preheated before
the oxygen, water, and slurry streams are passed to the SCWO system. The SCWO
system is operated at 250 bar to remain well above the critical pressure (221 bar), thereby
avoiding major fluctuations in properties. A low-sulfur Powder River Basin coal was
chosen for the model [9]. The coal exergy was determined by the methods outlined by
Kotas [10].
The regenerator/desalinator contains a set of staged heat exchangers, designed to best
transfer the enthalpy of the combustion products stream to streams where heating is
required. To understand its operation, it is best to follow both the schematic in Fig. 2 and
the temperature profiles in the various heat exchangers shown in Fig. 3. The products
enter at the top right of Fig. 2 and heat the desalinated water and coal slurry to
supercritical temperature in a regenerator, before they enter the SCWO system. The
products stream temperature profile inflects as it passes over the vapor dome. Therefore,
the desalination process is split into a two-pressure-level system in order to best match
the temperature profile of the products stream in the intermediate pressure (IP) heat
exchanger. Two pressure levels are also necessary to provide enough desalinated water
for the SCWO system without creating a pinch point, and to prevent salts from
precipitating in the brine if too much water is evaporated.
Figure 1: Diagram of a SCWO system combined with a heat engine
powered by the hot combustion products, an air separation unit, a
slurry preparation unit, and a combined regenerator/desalinator.
AQUIFER BRINEINJECTANT
Lift
Pump
Injection
Pump
COAL
Slurry
Prep.
AIRASU
NITROGEN
IP
Pump
LP
Pump
HP
Pumps
Slurry
Pump
Regenerator
IP Desal.
Products
Stream
To SCWO
System
Oxygen
LP Desal.
Q1
Q2
IP HX
LP HX
Brine Econ.
Slurry HX Slurry Regen.
Figure 2: Diagram of the regenerator/desalinator. Q1 and Q2 are transferred
to cooling water during startup and to the HRSG at steady state.
Figure 3: Temperature vs. enthalpy plot of the streams in the regenerator/
desalinator. (Line type denotes the streams in each heat exchanger; arrows
give the direction of each stream)
After the regenerator, the combustion products pass through the IP heat exchanger
before being mixed with the other outlet flows and re-injected. The IP heat exchanger
brings the IP brine stream under the vapor dome, evaporating some water while
concentrating the remaining brine. After flashing in the IP desalinator, the concentrated
brine is drained and used to preheat the low pressure (LP) brine in the LP heat exchanger.
This process brings the LP brine to the saturated-liquid state. That liquid is then
evaporated by the condensation process of the IP desalinator. The LP desalinator flashes
the partially evaporated LP stream, and again returns the concentrated brine to a heat
exchanger to preheat the coal slurry to near-critical temperatures. The condensation
process in the LP desalinator is used to preheat the oxygen stream to near-critical
temperatures. These two preheating processes are not completely matched in capacity to
cool the brine and condense the water vapor, so cooling water or some of the bypassed
aquifer brine is used to provide additional cooling as shown by Q1 and Q2.
The desalinated water streams are then over-pressurized to 450 bar to minimize the
effect of a pinch point in the regenerator from the slopes near the pseudocritical point and
create a better temperature match with the three streams in the regenerator. The pressures
in the desalination process were chosen at 110 bar (IP) and 100 bar (LP) to minimize
exergy destruction in the heat exchangers. The intermediate pressure of 110 bar was
chosen to match temperatures well with the pseudocritical point of the products. This is
about half the critical pressure, which results in a saturation temperature close to the
temperature of the products stream. The composition of the products stream has a lower
critical temperature than pure water, and lowers the temperature of the products stream as
it passes over the vapor dome. There is also a pinch point to consider as the aquifer brine
begins to change phase, chosen to exceed 25 K in temperature difference. With the
desalination pressures close to each other, this allows for a small temperature difference
in the IP desalinator condenser, where the temperature difference is constant as each
stream changes phase. This low temperature difference minimizes exergy destruction in
the process.
Air NitrogenOxygenLOX
Pump
HX 1
LP
HP IP
Turb.
HX 3
HX 3HX 2
High
Press.
Col.
Low
Press.
Col.
Q3
Q3
Figure 4: Diagram of the Air Separation Unit. HX 3 is shown broken up
with both streams in the right block cooled by the stream in the top block.
A LOX ASU was analyzed based on available designs, and further optimized for this
system [11, 12]. The architecture used is displayed in Fig. 4. The system includes three
pressure levels of air compression. One stream is pressurized to that of the high pressure
(HP) column, and cooled in a series of heat exchangers using the separated nitrogen and
oxygen streams. Another is over-pressurized, cooled, and then expanded across a throttle
for Joule-Thomson cooling. The final air stream is over-pressurized, cooled and then
sent through a turbine to drive the IP compressor and in doing so obtain additional
cooling. These streams are then sent throughout the system with a HP column containing
a condenser that acts as a reboiler for the LP column. Streams entering the LP column
are subcooled in heat exchanger (HX) 3 by the nitrogen product.
The SCWO system is modeled as a single combustion process because the outlet state
of the system only depends on the streams entering the system. However, one could
imagine a system similar to a gasifier, where the coal slurry enters a reformer with some
oxygen and supercritical water. The reformer breaks down the coal into a synthetic fuel,
the solids are separated, and the remaining oxygen and water are injected in a combustor
at near stoichiometric proportions (since both the fuel and oxygen are valuable
resources). The combustion products then transfer heat to a heat engine, since the
products need to remain pressurized for re-injection. Finally, the products are cooled in
the regenerator/desalinator heat exchangers and injected back into the saline aquifer.
Main Heat
Exchanger
Brayton Cycle
Rankine Cycle
Work
Work
Heat Rejection
from Condenser
COMBUSTION
PRODUCTS
TO REGEN/
DESAL
Figure 5: Heat engine layout for the plant, including a helium Brayton
cycle, bottomed by a Rankine cycle.
The heat engine, depicted in Fig. 5, consists of a helium Brayton cycle heated by the
combustion products in the main heat exchanger, and bottomed by a conventional (water-
based) Rankine cycle. Helium was chosen as the working fluid because of its high
thermal conductivity and specific heat ratio. Both are important to aid heat transfer with
the combustion products. The helium flow rate is chosen to match the capacity of the
products stream in the main heat exchanger. To reduce creep stresses at this high
temperature, the pressures of the streams entering the heat exchanger are matched. The
Rankine cycle pressure is specified to ensure that the quality of the water exiting the
turbine is greater than 0.88. Maintaining a pinch point difference in the HRSG of 17 K
dictates the water flow rate in the Rankine cycle.
There are a number of parameters that are variable within the heat engine and
regenerator/desalinator. These systems are highly coupled and therefore their parameters
are difficult to optimize to find the maximum plant efficiency. For example, the Brayton
pressure ratio determines the temperature of the helium entering the main heat exchanger
after compression. This temperature sets the total heat transferred, for a specific
effectiveness, in the main heat exchanger. Meanwhile, a pinch point is possible near the
low-temperature side of the main heat exchanger, due to the concavity in the products
stream temperature profile as it nears the pseudocritical point. In addition, the products
stream must have a high enough temperature to heat the slurry and water to supercritical
temperature in the regenerator. However, as the slurry and water temperature exiting the
regenerator increases, additional desalinated water is required to moderate the SCWO
system. This, in turn, raises the products stream heat transfer in the IP heat exchanger,
increasing the chance of a pinch point within the regenerator and IP heat exchanger.
These competing factors, among others, must be optimized to find the most efficient
operating conditions for the system.
Model Results
The LOX ASU was found to have a work requirement of 1.4 MJ/kg-O2, including
pressurization to the system combustion pressure of 250 bar. Figure 6 shows the oxygen
work intensity of the LOX ASU and a range for a state-of-the-art GOX ASUs with gas
compression to varying pressures. The gaseous work requirement was found from
published data at 1.7 bar [12] and was compressed further with a polytropic efficiency of
90%, in four intercooled stages. The number of intercooled stages could make the
gaseous system a more, or less, work-intensive option compared to the liquid-pumped
ASU. The oxygen leaving the ASU at full pressure from a LOX/pumped plant is near
ambient temperature, while a GOX stream could be at a much higher temperature from
the compression process, if not well intercooled. This hot oxygen could be advantageous
to reduce preheating loads or eliminate the need for preheating in the
regenerator/desalinator. However, the low heat capacity of the oxygen requires less than
14% of the heating load from the LP desalinator condenser as seen in Fig. 3. In addition,
decreased levels of oxygen preheating would coincide with increased oxygen work
intensity as intercooling is reduced. Therefore, the differences in the systems are quite
small, with both gaseous and liquid ASUs appearing as good options for this plant.
0 100 200 300 400 500
0.8
1
1.2
1.4
1.6
Exit Pressure, bar
Oxygen Inte
nsity, M
J/k
g O
2
Liquid Pumped ASU
Compressed Gas IPCC ASU-High
Compressed Gas IPCC ASU-Low
Figure 6: Plot of the oxygen intensity of the modeled liquid pumped ASU
versus the high and low range of published data [12] for a gaseous ASU
with intercooled compression. The dashed line is the nominal operating
pressure of the SCWO plant.
The temperature vs. products-specific enthalpy plot shown in Fig. 3 displays the
functioning of the regenerator/ desalinator. The IP desalinator and regenerator are
designed to minimize temperature differences while cooling the products stream
(decreasing irreversibility). The rest of the heat exchangers are heated by the cooling of
the concentrated brines and condensation of water vapor, shown underneath the IP
desalinator in Fig. 3, as described in the Model section. The flow direction of each
stream in the heat exchangers is shown by the arrows on the plot. A pinch point of 25 K
was found in the IP desalinator, and 11 K in the regenerator.
After an exergy analysis—investigating the irreversibility in each plant component —
it was found that large exergy destruction occurs in the LP desalinator and brine
economizer. This destruction is because the extra cooling loads Q1 and Q2, shown in
Fig. 2, remove enthalpy from a valuable resource at high temperature and provide no
useful service. Given that the LP desalinator saturation temperature is above that of the
water in the HRSG, it is possible to heat a portion of the HRSG evaporator while
condensing steam in the LP desalinator. The brine economizer can also transfer heat to
the HRSG economizer. This heat integration, shown in Fig. 7, may be difficult to
accomplish during startup procedures in a practical installation. Therefore, the LP
desalinator should be setup to be cooled by bypassed aquifer brine during startup
procedures, while using the heat integration during steady state operation for maximum
efficiency.
Main Heat
Exchanger
Brayton Cycle
Rankine Cycle
Work
Work
Heat Rejection
from Condenser
COMBUSTION
PRODUCTS
TO REGEN/
DESAL
Q1 Q2
Figure 7: Heat engine schematic of the plant at steady state conditions,
including heat input to the economizer and boiler from the
regenerator/desalinator as depicted by Q1 and Q2.
The power balance for a 500 MW plant is shown in Table 1. An overall efficiency of
36.7% was found. This efficiency includes all energy penalties for oxygen separation
and pressurization of the waste stream to the injection pressure.
Figure 8 shows the exergy distribution of the incoming coal for the SCWO plant
under startup and at steady state conditions. During startup, roughly 25% of the fuel
exergy leaves as net work out and the rest is destroyed in the various sections of the plant.
A large amount of exergy destruction occurs in the regenerator/desalinator since the heat
exchangers are inherently irreversible due to temperature differences between the
streams. After heat integration is incorporated, the exergy distribution, shown by the
green bars on the right in Fig. 8, display an 11% decrease in exergy destruction in the
regenerator/desalinator. Almost all of that exergy is transferred into increased work
output. The integrated system has an efficiency of 36.7% on an LHV basis. The brine
economizer is able to provide 30% of the heating in the economizer and 56% of the
heating in the evaporator. This extra heating permits a larger water mass flow rate in the
Rankine cycle, providing a higher net work output from the plant. In addition, a closer
temperature match exists with the helium in the superheater as a result.
Table 1: Work flow of power plant under steady state conditions.
Component Power (MW)
Brayton Cycle
Compressor -314.4
Turbine 527.1
Net 212.7
Rankine Cycle
Condensate Pump -0.03
Feed Pump -2.69
Turbine 461.6
Net 458.8
ASU -146.1
Water Pumps -25.5
Overall Plant 500.0
Firing Rate (LHV basis) 1364.1
Efficiency ( % LHV) 36.7%
Figure 8: Clustered bar plot comparing the plant under startup (left) and
steady state (right) conditions with heat integration, leading to increased
efficiency.
This SCWO plant is compared to other coal technologies with CO2 capture in Fig. 9.
The SCWO plant is a more efficient option than any of the three currently available
technologies using coal, including subcritical and supercritical PC with post-combustion
capture, and IGCC with pre-combustion capture [13]. These plants only capture 90% of
the CO2, and do not capture other toxins such as mercury. The SCWO plant captures
100% of the CO2 from the plant with zero emissions to the atmosphere, including trace
species. The SCWO injectant is also the most secure given that it is pre-equilibrated with
the aquifer brine.
Figure 9: Overall efficiency of current coal technologies with 90% CO2
capture and a SCWO plant with 100% capture and equilibrated.
Further increases in efficiency could follow from the insights gained from the exergy
analysis. Currently, the largest exergy destruction occurs due to irreversibilities during
combustion. The important variables that affect exergy destruction from combustion in a
SCWO system are the reactant and product temperatures. Figure 10 displays the
percentage of incoming fuel exergy destroyed by combustion against these two
parameters. This contour plot illustrates the Extreme States Principle as outlined by Teh
[14]. Teh found that combustion exergy destruction could be lowered by going to high
energy states, at high pressures and temperatures, before combustion. In the case of
SCWO systems, exergy destruction from combustion has a weak dependence on
pressure. This causes reactant and product temperatures to be the dominant factors.
Figure 10 shows the current system, marked by a circle on the plot, operating at a
maximum product temperature of 1600 K. Material constraints in the main heat
exchanger limit the product temperature in this case. The dashed arrow in Fig. 10 shows
that a reactant temperature increase of roughly 75 K at a fixed product temperature could
lead to a 5% reduction in exergy destruction, while another 200 K increase would be
necessary to achieve another 3% reduction. The heat integration effects reveal that most
of this reduction could translate into a similar increase in net efficiency. Future analysis
will attempt to achieve such changes while balancing the optimum operation of the plant
as outlined in the System Model section. Future research will also include a sensitivity
analysis for the important parameters of the system.
Conclusions
The modeled system has furthered understanding of the SCWO power plant initially
proposed under this project. Either a liquid-pumped or gaseous ASU was found to be
appropriate for the modeled SCWO plant. Heat integration was found to substantially
increase the overall efficiency of the plant. The efficiency of the plant was found to be
36.7 %, higher than current low-emission coal technologies. A path towards further
efficiency improvements through combustion exergy destruction was also identified.
Figure 10: Combustion exergy destruction (% of fuel exergy) as a function
of reactant and product temperature at 250 bar. The present system state is
marked with a circle. The x denotes the state that would lower exergy
destruction by 5% if the reactant temperature is increased by 75 K.
The analysis of the full SCWO system shows that it is a potential option for the
efficient use of coal in electricity generation with zero emissions to the atmosphere. The
SCWO technique has the added advantage of injecting a pre-equilibrated mixture of
aquifer brine and combustion products, including CO2. This mixture is denser than the
original aquifer brine, allowing for a more secure storage strategy than neat CO2
injection. This type of injection may allow for decreased monitoring and liability costs
for CO2 storage in saline aquifers. It may also lead to an increased number of aquifers
available for CCS, if it is found that a suitable caprock is not required for effective long-
term storage. Increasing the potential number of storage aquifers suitable for CO2 storage
could reduce costs of establishing that an aquifer will retain the majority of the injected
CO2, lowering costs associated with permitting. Injection sites could also be located
closer to CO2 sources, reducing transportation costs. Reducing the transmission, storage,
and monitoring costs all could have significant impacts on reducing greenhouse gas
emissions, however, the largest cost of CO2 abatement is from the capture process. This
technology could have a large impact on the cost of CO2 capture, since separation
processes are minimized as well as their associated energy penalties, resulting in an
efficient power plant.
Publications
1. Heberle, J.R., & Edwards, C. (2009). Coal energy conversion with carbon
sequestration via combustion in supercritical saline aquifer water. International
Journal of Greenhouse Gas Control, 3 (5), 568-576.
2. Mobley, P.D., Pass, R.Z., Edwards, C.F. Exergy Analysis of Coal Energy
Conversion With Carbon Sequestration via Combustion in Supercritical Saline
Aquifer Water. 2011 Energy Sustainability Conference. Accepted for publication.
Contacts (Systems Analysis)
Christopher F. Edwards: [email protected]
Paul D. Mobley: [email protected]
Rebecca Z. Pass: [email protected]
References (Systems Analysis)
1. EPRI, 2007. The Power to Reduce CO2 Emissions.
2. Benson, S. C. (2005). Underground geological storage. In B. D. Metz, IPCC
Special Report on Carbon Dioxide Capture and Storage (pp. 195-276).
Cambridge: Cambridge University Press.
3. Burton, M. B. (2007). Eliminating buoyant migration of sequestered CO2 through
surface dissolution: implementation costs and technical challenges. Society of
Petroleum Engineers Annual Technical Conference and Exhibition. Anaheim, CA,
USA.
4. Wellig, B., 2003. Transpiring-wall reactor for supercritical water oxidation.
Doctoral thesis no. 15,038, Swiss Federal Institute of Technology (ETH) Zurich.
Available http://www.e-collection.ethz.ch.
5. Prikopsky, K., Wellig, B., & von Rohr, R. (2007). SCWO of salt containing
artificial wastewater using a transpiring-wall reactor: Experimental results.
Journal of Supercritical Fluids , 40, 246-257.
6. Donatini, F. (2009). Supercritical water oxidation of coal in power plants with
low CO2 emissions. Energy , 34, 2144-2150.
7. Armellini, F. T. (1994). Precipitation of sodium chloride and sodium sulfate in
water from sub- to supercritical conditions: 150 to 500 C, 100 to 300 bar. The
Journal of Supercritical Fluids , 7 (3), 147-158.
8. Hangx, S. (2005). Behavior of the CO2-H2O System and Preliminary
Mineralisation Model and Experiments. Utrecht University, Department of Earth
Sciences.
9. Argonne Premium Coal Sample. (n.d.). Retrieved August 10, 2010, from Argonne
National Laboratory: http://www.anl.gov/PCS/report/part2.html
10. Kotas, J. (1985). The Exergy Method of Thermal Plant Analysis.
11. Howard, H. (2008). Patent No. 7,437,890. US.
12. Thambimuthu, K. S. (2005). Capture of CO2. In B. Metz, IPCC Special Report on
Carbon Dioxide Capture and Storage (pp. 105-178). Cambridge: Cambridge
University Press.
13. DOE/NETL. (2010). Cost and Performance Baseline for Fossil Energy Plants,
Volume 1: Bituminous Coal and Natural Gas to Electricity.
14. Teh, K.-Y. (2007). Thermodynamics of Efficient, Simple-Cycle Combustion
Engines. PhD Thesis, Stanford University, Palo Alto, CA.
Area 2: Supercritical coal reforming (Mitchell)
The proximate and ultimate analyses of Wyodak coal, the sub-bituminous coal from
Wyoming selected for study, are presented in Table 1. On a dry basis, the coal has a
lower heating value of 28.2 MJ/kg. Samples of the coal were ground and size classified
to obtain particles in the size range 75 – 106 m for use in the SCW reforming tests.
Table 2: Proximate and Ultimate Analyses of Wyodak Coal
Proximate Analysis Ultimate Analysis
Property Wt-% Property Wt-%
Fixed carbon 35.06 Carbon 51.42
Volatile matter 33.06 Hydrogen 4.16*
Moisture 26.30 Nitrogen 0.69
Ash 5.58 Sulfur 0.32
Chlorine < 0.01
Oxygen 11.53#,*
Ash 5.58
*excludes moisture #determined by difference
When coal particles are heated, volatile matter (organic coal decomposition products)
is released and char particles are produced. The volatile matter consists of evolved gases
(CH4, CO, CO2, H2, H2O, and light hydrocarbons (C2H2, C2H4, etc.)), and tars. The coal
decomposition products classified as tars are organic compounds having molecular
weights greater than benzene (78 kg/kmol) and which condense at low temperatures. The
tars released from the coal (the primary tars) are oxygenated compounds. These are
converted to secondary tars (phenolic compounds and olefins) and tertiary tars
(polycyclic aromatic hydrocarbons, PAH) in steam gasification environments. The
secondary tars are also converted to PAH in such environments. In SCW environments,
the tars and PAH hydrolyze rapidly, yielding CO, CO2, H2, and H2O, primarily, with
smaller amounts of CH4 and other light hydrocarbons. These species are all completely
miscible in supercritical water.
The char particles remaining subsequent to coal devolatilization react with water in
SCW environments at a rate somewhat slower than the rate that the tars are converted to
dissolved gases. Consequently, the size of the reformation reactor is governed by the
conversion time of the coal char in SCW environments. Our experimental efforts are
devoted towards providing this information.
Autothermal Reformer Operation
Calculations using the thermodynamic model developed in Area 1 activities indicate
that the residual sensible energy in the stream leaving the heat exchanger where energy is
transferred to the heat engine can be used to preheat the water fed to the reformer to a
temperature as high as 700 K. By preheating the reactants for gasification, the amount of
oxygen required for autothermal operation of the reformer is reduced and the heating
value of the synthesis gas is increased. Shown in Fig. 11 is the required amount of
oxygen per 100 kg of coal to operate an autothermal reformer at 647.3 K as a function of
total pressure when the water is preheated to 700 K. The required amount of oxygen is
relatively insensitive to the total pressure for a given solids loading (the coal weight
fraction in the slurry).
At any selected reformer operating pressure, as the solids loading is decreased, more
slurry water needs to be heated. With preheated water, for a specified reformer operating
pressure, the amount of oxygen required to keep the gasification temperature at 647.3 K
by partial oxidation of the coal decreases as the solids loading decreases.
By preheating the water supplied to the reformer, less energy needs to be produced
from combustion of coal and volatilized species; a higher concentration of combustible
species exists in the synthesis gas compared to the no-preheat case. Also, the lower the
coal fraction in the slurry, the less oxygen required for autothermal gasification.
Consequently with preheated water, the lower the solids loading the higher the heating
value of the synthesis gas. This is demonstrated in Fig. 12, in which are shown
calculated values of the water partial pressure in the reformer as a function of the total
reformer pressure and solids loading. The ratio of the heating value of the synthesis gas
to the heating value of the coal (the HV ratio) is indicated for each solids loading curve.
220 240 260 280 300 3201.6
1.8
2
2.2
2.4
2.6
2.8
3
Ptotal
(atm)
O2 a
dd
ed
(km
ol/100kg
of
co
al)
Solids Loading = 0.150
Solids Loading = 0.200
Solids Loading = 0.300
Solids Loading = 0.400
Figure 11: Required oxygen for autothermal operation of the
reformer at 647.3K with preheating of the reformer feed water.
220 240 260 280 300 32050
100
150
200
250
300
Ptotal
(atm)
PH
2O (
atm
)
Solids Loading = 0.150
Solids Loading = 0.200
Solids Loading = 0.300
Solids Loading = 0.400
(Syngas HV)/(Coal HV) = HV Ratio = 0.97
HV Ratio = 0.92
HV Ratio = 0.87
HV Ratio = 0.84
Figure 12: Heating values of synthesis fuel at 647.3K
with water preheating.
When evaluating the energy requirements of the reformer, real gas properties were
used for all species for which data are available (H2O, O2, N2, CO2, CH4) when
determining the equilibrium composition of the synthesis gas. For species for which real
gas data are not available (H2S, HCl, Cl2, CO, COS, C2H2, C2H4, C2H6, NH3, NO, NO2,
SO2) the Peng-Robinson equation of state was used to describe their P-v-T behavior.
Shown in Fig. 13 is a comparison between predicted synthesis gas compositions
assuming ideal gas behavior and real gas behavior at 290 atm and 647.3 K, a supercritical
water environment. The elemental composition of Wyodak coal, shown in Table 2, was
used to determine the atom populations. The scale of the vertical axis is logarithmic.
Noticeable differences are significant. In particular, when real gas behavior is taken into
account, the predicted synthesis gas has less CO and H2 and more CH4 than the synthesis
gas predicted assuming ideal gas behavior.
Figure 13: Predicted equilibrium evolved gas composition at 290 atm, 647.3 K.
Atom populations were determined from the elemental analysis for Wyodak coal,
presented in Table 2.
Coal Char Reforming
No gasification model has been developed that permits the accurate prediction of
particle mass loss rates under supercritical water conditions. Based on previous work, we
have developed the heterogeneous reaction mechanism presented in Table 3. In the
mechanism, water dissociatively chemisorbs onto free sites (Cf) on the carbonaceous
surfaces yielding adsorbed OH and H (C(OH) and C(H), respectively). Adsorbed O
(C(O)) is also formed as C(OH) dissociates on the surface. The adsorbed species are
involved in surface reactions that result in the release of H2 and CO, primarily, with
smaller amounts of CH2, CH and COH (which rearranges to HCO) also released. Once
in the gas phase, the CH and CH2 are involved in homogeneous reactions that result in
the formation of methane. In order to accurately account for the changes in the gas phase
composition due to chemical reaction, a detailed homogeneous reaction mechanism is
required. In order to avoid this complexity, we assume that the global reactions shown in
Table 4 are kinetically limited, and that the H2/O2 system of reactions is equilibrated.
In our approach, the Arrhenius parameters for the reaction rate coefficients of the
heterogeneous reactions in Tables 3 and for the rate-limiting homogeneous reactions in
Table 4 are determined by adjusting the parameters to provide agreement with data
obtained in experiments performed in this project. The rate coefficients for the reverse
reactions of the rate-limiting homogeneous reactions are determined from the equilibrium
constants and the forward rate coefficients. Equilibrium constants for the homogeneous
reactions are determined from data provided in the JANAF Thermochemical Tables and
fugacity coefficients to account for non-ideal gas behavior.
The intrinsic chemical reactivity of the carbonaceous material is defined as the rate
that a carbon atom is removed from the bulk. Based on the heterogeneous reaction
mechanism shown in Table 3, carbon reactivity (Ri,C, in kg/m2-s) is given by
RiC k4 C(H) 2 k6 C(O) k7 C(H) k8 C(OH)
where the brackets denote the concentration of the species enclosed.
Table 3: Reduced heterogeneous reaction mechanism for char reactivity to H2O
2Cf + H2O <==> C(OH) + C(H) (1)
C(OH) + Cf <==> C(O) + C(H) (2)
C(H) + C(H) <==> H2 + 2Cf (3)
C(H) + C(H) ==> CH2 + Cf (4)
C(H) + C(OH) ==> H2 + C(O) + Cf (5)
C(O) + Cb ==> CO + Cf (6)
C(H) + Cb ==> CH + Cf (7)
C(OH) + Cb ==> COH + Cf (8)
Table 4: Rate-limiting global and equilibrium homogenous reactions
involved in methane formation and HCO destruction
Rate-limiting global reactions
CH + 2 H2O <==> CH4 + O + OH (9)
CH2 + H2O <==> CH4 + O (10)
CHO + H2O <==> CO2 + H2 + H (11)
Equilibrated reactions
H + OH + M <==> H2O + M (12)
H + H + M <==> H2 + M (13)
O + O + M <==> O2 + M (14)
If the gas phase species reach a state of chemical equilibrium, the above mechanism
will predict the expected behavior at both low and high pressures. Shown in Fig. 14 are
predicted synthesis gas compositions for a gasifier operated under autothermal, SCW
conditions at 290 atm and 647.3 K, with and without preheat. Atom populations were
based on the composition of Wyodak coal. Compared to the synthesis gas produced in a
gasifier operated under typical autothermal, Integrated Gasification Combined Cycle
(IGCC) gasification conditions (29 atm, 1700 K yielding primarily CO and H2), the
synthesis gas of the SCW gasifier has a considerably higher methane fraction with little
CO and H2.
Figure 14: Predicted equilibrium evolved gas compositions. SCW gasification at
290 atm, 647.3 K. IGCC gasification at 29 atm, 1700 K. Atom populations were
determined from the elemental analysis for Wyodak coal, presented in Table 2.
Experimental Facility
An experimental facility was designed, built and assembled that permits the
acquisition of data needed to determine the rate coefficients for reactions in Tables 3 and
4 that are suitable for supercritical water conditions (P > 218 atm and T > 647K). The
focus point of the facility is an entrained flow reactor that can be pressurized up to 340
atm at temperatures up to 1000 K. The reactor is 15 meters long, sufficient for residence
times up to 15 minutes when the coal-water slurry feed contains 20% solids and the total
mass flow rate (slurry plus supercritical water (SCW) flow rates) is 20 g/min. The
residence time in the reactor is varied by controlling the solids content of the slurry and
the flow rate of the water supplied. The feed water is pressurized and preheated to the
test conditions before the solids slurry and any oxygen are admitted. Oxygen is added so
that the heat release from partial oxidation of the solids can supply energy for
autothermal gasification.
The stream leaving the flow reactor is quenched and dumped into a water bath where
the solids are collected and analyzed to determine the extent of conversion. A side-
stream is directed to a gas analyzer, purchased from Rosemount Analytic, which
measures in real time the concentrations of four species of interest: CO, CO2, CH4, and
H2. The side-stream can also be directed to the sampling loop of our gas chromatograph
to permit the measurement of the concentrations of such other species as N2, O2, H2S, and
SCW (no preheating) SCW (preheating)
NH3. A schematic of the experimental facility is shown below in Fig. 14. A picture is
shown in Fig, 15.
Figure 14: Experimental setup for coal gasification under supercritical conditions.
A water pump, made by SC Hydraulic, is used to pressurize and supply water to the
system. This pump is driven by air (at 120 psi (8.17 atm)) supplied from a high-pressure
air tank. The air to drive this pump is regulated by a FESTO air controller to control the
water pressure. The water supplied to the reactor is preheated using three Chromalox
MaxiZone cartridge heaters surrounded by aluminum powder and calcium silicate
insulation. The water pump that supplies water to the slurry is manufactured by Haskel,
and the pump-driving air is controlled by another FESTO air controller.
A high-pressure piston-accumulator purchased from High Pressure Equipment
Company is used to deliver the coal-water slurry to the reactor. The slurry is loaded on
one side of the piston-accumulator by a peristaltic pump before the start of an
experiment. Pure high-pressure water from the Haskel pump delivers the loaded slurry
to the reactor by moving the piston from the pure water side toward the slurry side of the
accumulator. For slurry loadings at low pressures, a Vector peristaltic slurry pump is
used instead of the Haskel pump.
The water pumps only have a small volume per stroke compared to the overall
amount of water required for experiments consequently, the pumps need to stroke
frequently to supply the required flow rates of pressurized water. The fluctuations in the
flow due to these pump strokes are dampened by using pulse dampening accumulators
between the water pumps and the pressure transducers. The accumulators contain a
nitrogen-filled bladder inside a steel housing. The water pressure fluctuations are
dampened by expanding and contracting the bladders. The bladder accumulators were
purchased from HyDac.
Figure 15: A picture of the experimental facility.
The oxygen that is supplied to the reactor is pressurized using a SC Hydraulic gas
booster that is driven by air from a high-pressure air tank at 150 psi (10.2 atm). The
pressurized oxygen is stored in a stainless steel cylinder and flows through a high-
pressure oxygen flow controller made by Bronkhorst High-Tech. The preheated pure
water and the slurry from the piston accumulator are mixed with high-pressure oxygen at
a mixing junction, located at the starting point of the coiled reactor.
The coil-shaped 15-meter long reactor (see Fig. 16) is made of Inconel-625 to endure
high-pressure, high-temperature environments. The reactor is submerged in a bath filled
with sand (pulverized aluminum oxide) and sixteen Watlow Firerod cartridge heaters.
The heaters are used to heat the air used to circulate the sand particles in order to
maintain a uniform sand bath temperature. These heaters are capable of maintaining the
sand bath at temperatures up to 1000 K. A two-channel heating system control station
was developed to control temperatures of the preheater and the reactor sand bath.
The flow leaving the flow reactor is rapidly quenched in a cool-down tube in order to
prevent further reactions. To decrease the cool-down time, the tube diameter at the
reactor exit is decreased in order to increase the velocity of the flow and improve the
overall heat transfer by convection. Cooling is provided by submerging the cool-down
tube in a chamber filled with flowing cold water. Calculations indicate that the 1-meter
long cool-down tube is sufficient to freeze all chemical reactions in less than 20 seconds.
With this arrangement, the coal reaction time is essentially the residence time of the
particles in the 15 meter coiled reactor.
Figure 16: The 15-m coiled tubular reactor.
The flow contains liquid, gas products and solid particles, which must be separated.
The gas is separated from the liquids and solids in the separation chamber, a high-
pressure cylinder made by High Pressure Equipment. At this stage, the flow is still at a
high-pressure but it is at a low temperature after being cooled in the cooling chamber.
Due to the limited volume of the separation chamber, during warm-up the flow bypasses
the separator. The flow is directed through the separator only during actual tests. High-
pressure HIPCO valves are used to direct the flow.
The entire system pressure is controlled by a Tescom back-pressure-regulator. A
Marsh Bellofram air controller supplies sufficient air to the back-pressure-regulator to
maintain system pressures up to 374 atm.
The gas leaving the separation chamber is directed to an analyzer where the gas
composition is measured. Gas analysis is made using an Emerson Process Management
X-STREAM gas analyzer that has four channels: CO (0~40%), CO2 (0~40%), CH4
(0~20%), and H2 (0~50%). For measurements of gaseous products other than these four
products, our Varian gas chromatograph is used.
Air supplies to drive the water pumps, the oxygen booster, and the HIPCO valves are
controlled by ASCO solenoid valves, which are operated by OPTO22 modules on a data
acquisition box. The data acquisition/system control box consists of a data acquisition
board (DaqScan 2001) combined with extension boards (DBK 207/CJC, DBK 208 and
DBK 11A) for analog inputs/outputs and digital inputs/outputs from IOtech. Signals
from the Bronkhorst oxygen flow controller, the FESTO air controllers, and
thermocouples are also acquired on these boards. All inputs and outputs are managed
remotely from a dedicated computer. The automated system control and data
measurement programs are written in LabVIEW (National Instruments).
Continuing Research (Supercritical water reforming)
Coal conversion tests in supercritical water environments are underway. Extents of
coal conversion are being measured at selected residence times in various SCW
environments. The compositions of the synthesis gases produced are being measured as
well for each of the environments. The data will be used to determine rate coefficients
for the reactions in Tables 3 and 4. In order to help define the rates of some of the
reaction pathways, reactivity tests are also being performed in our pressurized
thermogravimetric analyzer.
In March 2010, the high-pressure compressor that supplies the high pressure air to the
pumps and solenoid valves broke, preventing the SCW facility to be operated.
Diagnosing the reason for the failure and deciding whether or not the compressor was
repairable, securing funds to either repair or replace the compressor, purchasing a new
compressor and installing it took a considerable amount of time. It was not until
February 2011 that the SCW facility was operational, permitting experiments to continue.
References (Supercritical water reforming)
1. Churchill, S.W., and M. Bernstein, J. Heat Transfer, 99, 300, 1977.
Contacts (Supercritical water reforming)
1. Reginald E. Mitchell: [email protected]
2. BumJick Kim: [email protected]
Area 3: Synthesis Fluid Oxidation and Heat Extraction (Edwards)
Supercritical Combustor Development
The next task of the project is the development of a supercritical water combustor.
Our goal is to demonstrate stable combustion in supercritical water with near-
stoichiometric amounts of fuel and oxidizer and with high volumetric firing rate. These
conditions are desirable for a combustor integrated into the plant described above.
Design and Layout of Experimental Apparatus
Wellig has identified six tasks that must be performed by a SCWO experimental
apparatus: reactant storage, pressurization, preheating, reaction, pressure let down, and
effluent discharge [1]. Due to the firing rates we intend to use and the amount of fluid we
will preheat, cooling of effluent is also a concern, since otherwise the ambient
temperature in the laboratory could rise to an unacceptable level.
Fig. 17 is a simplified schematic showing the major hardware for the supercritical
combustor experiments. At left are components to store and deliver pressurized fuel,
oxygen, and water. In the center there is an electric resistance heater ahead of the reactor
vessel. Combustion occurs above the heater in the lower section of the high-pressure
stack. The upper section of the stack is a dilution cooler, where ambient temperature
water is introduced to begin cooling the combustor effluent. At right is a condenser to
complete cooling of products and reject thermal energy from the lab. Each subsystem is
discussed in detail below. Several changes have been made in the configuration of these
systems since the last annual report.
Figure 17: Diagram of key components of the supercritical water combustor
experiment. Reactant storage and pressurization components are shown to the left
of the vertically oriented high pressure stack. At right are components to remove
waste heat from the lab and separate the effluent components for disposal.
Figure 18: Liquid fuel delivery system.
The fuel system is shown in Fig. 18. Prior to running the combustor, a low-pressure,
air-powered pump (not shown) transfers fuel into a bladder accumulator in the laboratory.
This accumulator was sized to hold enough fuel for a four-hour run at a maximum firing
rate of 50 kW. During operation, the gas side of the bladder is pressurized with nitrogen
to feed fuel into the combustor. This strategy avoids the use of a high-pressure fuel
pump. The nitrogen is supplied from a six-pack of gas cylinders, pressurized with a two-
stage, air-powered booster, and regulated to a pressure above the combustor operating
pressure. Most of the (known) pressure drop from the accumulator pressure to the
combustor pressure occurs across a metering valve. This needle valve is operated
remotely with a stepper motor, so that fine changes in the valve’s flow coefficient allow
for adjustment of the fuel flow rate. The fuel line also has a normally-closed solenoid
valve for positive fuel shut-off. The fuel then flows through a spur-gear flow meter and a
check valve. Next, the fuel mixes with an adjustable amount of water so that the fuel
mass fraction entering the combustor can be varied.
Figure 19: Oxygen delivery system.
Figure 19 shows the oxygen system. Oxygen is fed from a six-pack of cylinders in
the oxidizer galley. Like the fuel bladder, this quantity of oxygen was chosen since it has
capacity for a four-hour run. The oxygen is regulated to 600 psi and flows into the lab,
where it is pressurized by a two-stage booster similar to the nitrogen booster in the fuel
system. To provide a steady flow of gas, the output of the booster is connected to a
plenum and a regulator. By pressurizing this plenum above delivery pressure, the
regulator just after it can deliver oxygen at an approximately constant pressure while the
piston in the booster cycles. Next are valves for manually isolating and bleeding
different parts of the system. A stepper-driven metering valve is used for remote control
of the flow rate, similar to the fuel system. The tube downstream of these valves is fitted
with a water cooling jacket. The jacket is used to reduce temperature variations in the
oxygen as the plenum blows down, facilitating repeatable metering and flow
measurement downstream. The remaining components are similar in function to those in
the fuel system. There is a remote solenoid valve for positive oxygen shut-off. The flow
meter consists of a needle valve in parallel with a differential pressure transducer. This
arrangement is similar in operation to a laminar flow element. The needle valve is set to
create an appropriate pressure drop, but is not adjusted during the experiment. A
thermocouple is located just upstream of the flow meter so that the latter’s readings may
be corrected for temperature variations. A gauge pressure transducer after the flow meter
arrangement serves a similar purpose. Finally, there is a check valve and a mixing tee.
While water will almost always be mixed with the fuel, initial combustion experiments
will involve injection of neat oxygen. The tee gives the option of mixing water with the
oxygen later, which is a way to affect mixing in the combustor by changing the relative
density and viscosity of the fuel/water and oxygen/water streams.
The low-pressure water system can be seen in Fig. 20. At the top left of the diagram,
water enters from a domestic cold water supply and proceeds through a 20 micron
particle filter. The line goes to a manual three-way valve that accepts either fresh water
from the supply or recycled water from the drain tank. This valve directs the water to a
rotameter (shown as FI for flow indicator) to ensure that the flow rate through the
deionizer is not more than 2 gpm. This flow rate is the recommended limit for effective
operation of the deionizer. Deionization is used to reduce corrosion in the heaters and
high-pressure stack. After the water is treated in the deionizer, it enters a 400 gallon
supply tank. This tank is filled before each set of combustion experiments and holds
enough water for a four-hour experiment.
The supply tank contains a submersible pump that sends water up into the head tank
with a flow rate higher than is necessary for the experiment. The excess water overflows
back into the supply tank, while water for the experiment is gravity fed through a 20
micron particle filter and isolation valve before entering a triplex plunger pump rated for
5000 psi (345 bar). This positive-displacement pump provides all of the high-pressure
water for the system. From the pump, the water goes to the high-pressure water manifold
shown in Fig. 21. Much like the head tank, the water manifold is supplied by the triplex
plunger pump with more water than is necessary for the experiment. Excess water
returns to the supply tank from the manifold back pressure regulator (BPR) described
later. There is a liquid level switch in the supply tank that controls the submersible pump
to keep it from running dry as the water level drops in the supply tank. The head tank
also has a liquid-level switch that controls the motor driving the triplex plunger pump to
prevent it from running dry.
The outlet stream from the high-pressure apparatus re-enters the low pressure system
near the top left of Fig. 20 after depressurization in the stack BPR. The main components
of the fluid here will range from water at ambient temperature during start-up and shut-
down to a mixture of saturated steam (maximum vapor fraction of 0.5) and carbon
dioxide during combustion. The flow will pass through the main condenser to cool the
stream to near ambient temperature. The main and residual condensers will use process
cooling water (PCW) from the building at an average of about 16°C and 30 gpm. The
CO2 and water vapor pass into the residual condenser to cool them below room
temperature. This prevents water vapor from condensing in the exhaust line running to
the roof. There is also a condensate drain in the vent line in case there is any
condensation due, for example, to cold outdoor temperatures. Thermocouples in both the
gas and liquid lines exiting the condensers are used to make sure that the condensers are
operating correctly. The lines connecting the two condensers are arranged to form a p-
trap and maintain a high water level in the main condenser for efficient heat transfer. The
bottom outlets from the two condensers send cooled water to the drain tank. This tank
has a centrifugal pump connected to its outlet with a liquid-level switch to make sure that
Figure 20: Low-pressure water handling system.
it does not run dry. The outlet of the pump passes through a check valve and a manual
three-way valve to send the water out the lab drain, or back through the deionizer and
into the supply tank to be used again.
The system for handling high-pressure water is shown in Fig. 21. These components
take filtered water from the low-pressure supply system, pressurize it, and control and
measure the flow rates of the several water streams that enter the high-pressure stack. At
the outlet of the triplex pump described above is a relief valve set to 5000 psi. After the
pump, the water passes by a bladder accumulator and through a manual needle valve.
Here an accumulator is employed in its usual role for pulsation dampening only, rather
than as both a reservoir and dampener as in the fuel system. The accumulator and valve
together act as a low-pass RC filter to attenuate high-frequency pressure waves. The cut-
off of the filter can be tuned by adjusting the valve. The water then enters a manifold,
where it branches out into seven lines. The manifold has a globe valve for each outlet so
that any lines that are not needed may be isolated. Excess flow from the pump leaves
through the manifold back pressure regulator (BPR) and returns to the supply tank. The
BPR is operated by an embedded PID controller that tracks the manifold pressure with a
transducer. This setup allows the BPR to keep a constant, set pressure in the manifold
regardless of changes in flow rate through each output line. This configuration decouples
the flow control in the output lines from each other.
Figure 21: High-pressure water handling system.
After the manifold, the water lines each have a stepper-driven needle valve, spur-gear
flow meter (except one), and check valve. Unlike the fuel and oxygen systems, each
water line does not have a solenoid shut-off valve. These are not needed since a small
amount of water flow past a mostly closed metering valve is not a problem in any of the
six lines. Furthermore, the needle valve selected for most of the lines is rated for
repeated shut-off by the manufacturer. After the check valves, two of the lines reach the
mixing tees described earlier for fuel and oxygen. The fuel and oxygen solutions and
three more water lines flow through the heaters and to the combustor. The last two lines
are for dilution cooling water. One goes directly to the cooling section of the high-
pressure stack. The other joins the line from the stack before the BPR.
There is a single outlet from the stack at the end of the dilution cooling section. It
runs to a second BPR. This BPR is responsible for holding the desired operating pressure
inside the combustor. Just ahead of it is a pressure transducer for feedback to the BPR
controller and a thermocouple to ensure that the fluid is below the rated maximum
temperature of the BPR. (This temperature limit is discussed below.) The manual three-
way valve after the BPR is for calibrating the water flow sensors. With the heaters off
and the stack cold, cold water is run through the pressurized system and out through this
valve. By measuring the mass of water output in a known time at several rates, the true
mass flow rate is correlated with the sensor output. The response is quite linear even at
low flow rates, despite the fact that this type of flow meter is intended for use with fluids
more viscous than water.
The BPR downstream of the combustor was initially rated to 650°F (617 K) by the
manufacturer. This is the highest temperature rated model produced by Tescom. The
difficulty in operating this assembly at high temperatures lies with the seals. After two
identical failures of a spring-energized, fluorocarbon radial c-seal, the failing seal was
replaced with an EPDM o-ring. This material is only rated to 250°F (394 K). To meet
this lower maximum inlet temperature, more dilution cooling was needed. With a
reasonable pressure drop from the water manifold to the stack BPR, the original 1/4 in.
dilution cooling line was not sufficient, even with the metering valve fully open. A new
1/2 in. line was installed to satisfy the increased demand. This new line is the right-most
water line leading from the manifold in Fig. 21. Presently it does not have a spur gear
flow meter as do the other lines. The flow rate is simply adjusted up or down depending
on the temperature of the mixed stream flowing into the BPR.
The heaters, shown assembled in Fig. 22, were designed with operational flexibility
in mind. They must heat at least five separate streams above the critical temperature of
water. A computer model was constructed to predict the necessary length of each line to
heat its stream to 700 K. The model operates under a constant wall temperature
assumption. This assumption is justified by the use of cast aluminum as the medium
between the electric heaters and tubes to be heated. The aluminum has a high thermal
conductivity and low spatial temperature variation has been observed in a similar heater
unit in use for another project in the research group. The model was modified for the use
of variable properties and assumes that the fuel/water line is entirely water. However, the
heat transfer model used could only give a rough estimate because it cannot account for
pseudoboiling in the tubes. Pseudoboiling occurs when the tube wall temperature is
above the pseudocritical point (the temperature of maximum specific heat at a given
pressure above the critical pressure) while the bulk of the flow is below the pseudocritical
temperature. At this condition the flow has a low-density supercritical boundary layer
and a high-density liquid core. Therefore, this regime operates very much like film
boiling at subcritical pressures. The supercritical layer does not conduct heat as well into
the core of flow, so deteriorated heat transfer occurs. The computer model was run twice
with different properties to give best- and worst-case scenarios. To work between these
two bounds, it was decided to construct the heaters with a large number of relatively short
tube passes. This design allows the total passage length for each of the five flows to be
varied between experimental runs by connecting together different numbers of tubes in
series.
The total power output of the heaters was chosen based on the enthalpy required to
preheat enough water, oxygen, and fuel to fire the combustor at rates up to 50 kW and
with mixed mean outlet temperatures from 1200 to 1800 K. About 40 ft. of electric
resistance cartridges of the chosen model are required to provide the desired 72 kW of
heating. For ease of fabrication, transport, and use, it was decided to use a larger number
of shorter cartridges. This choice works well with the choice of many short fluid tubes.
Due to the large numbers of heaters and tubes, two identical heater units were designed
and constructed. This geometry provides even heat transfer and convenient fluid tube
connections. Each unit is independently controlled by an embedded process controller.
This arrangement allows for the possibility of operating streams at different temperatures
to simulate various operating conditions in later experiments.
Figure 22: Photograph of the assembled heater units. Aluminum filler is
visible through the casting holes in the shells. At the end of each unit,
six cartridge heating elements and their wires are visible, along with
inner and outer circles of Inconel 625 fluid tubes.
Each unit is comprised of a steel shell with machined header plates welded to each
end. The end plates have holes for six 6 kW electric cartridges, 40 in. long, and twelve
¼‖ OD Inconel 625 fluid tubes. The heaters are inserted into thin-walled steel sleeves.
After insertion of the tubes and heaters, the shells were filled with cast aluminum. The
aluminum transfers heat from the heater cartridges to the fluid tubes. The tubes are 60 in.
long—20 in. longer than the heaters—to allow for tubes to be shortened when worn
connections need to be cut and replaced. (The tubes are not replaceable since they are
cast in the aluminum.) The tube ends were offset axially so they could be placed close
together without interference between adjacent fittings.
Calcium silicate was selected as the insulation for the heaters. Two-inch thick
insulation wraps around the heaters along with 2-in. thick end caps drilled for the tubes to
pass through. Any gaps in the hard insulation were filled with fiberglass insulation. The
heat loss to the environment from the heaters is less than 2%, based on a finite element
model of the heaters using a free convection heat transfer coefficient at the outer
insulation boundary.
A cut-away view of the high-pressure stack is shown in Fig. 23. The stack has three
major sections: the combustor, an elbow, and the dilution cooler. These sections are
connected with Grayloc-style flanges (in blue), which eliminates the need for welding
traditional flanges to the 2 in. thick vessel walls. They also allow the two flanges to be
placed close together without interference. The body of each section is machined from
solid Inconel 625 forgings. All fluid, thermocouple, and pressure fittings are cone-and-
thread high-pressure fittings to ensure proper sealing at elevated temperatures.
The combustor (the lower, vertical section in Fig. 23) has two functional zones. The
lower portion is the header, with inlet ports for the fuel/water solution, oxygen, and
water. These streams flow up through three coaxial passages formed by the outer vessel
and two nozzles. Nozzles of varying diameter may be used to change the relationships
between mass flow rates, fluid velocities, and Reynolds numbers at the burner. Swirlers
may also be inserted to affect flame stabilization. Near the center of the combustor are
two opposed sapphire windows for observation of the flame. (The viewing axis is normal
to the page, so only the inside of the far window is visible in this section.) In this
drawing the nozzles end just above the bottom of the window so that flame attachment at
the burner can be observed. Shorter nozzles could be substituted to observe flame
lengths longer than the window aperture. Above the windows there is a liner (shown in a
darker blue-gray than the main body). This perforated metal liner functions like that in a
gas turbine combustor, allowing relatively cool fluid (supercritical water in this case) to
enter from the outer annular passage, shielding the liner and the pressure wall from the
hot combustion products in the core flow. The pattern of 3/64 in. holes for this purpose is
shown. This liner extends beyond the combustor, through the flanged joint and into the
elbow.
The elbow section turns the flow and holds a third sapphire window to provide an
axial view of the flame in the combustor below. The window design (visible in this
view) is identical to that of the two windows in the combustor. The bolted flange holding
the window allows the window seals to be preloaded so they maintain a seal at the target
operating conditions of 700 K and 250 bar. There is an inlet near the window for
injection of 700 K water to prevent impingement of very hot (1800 K or more)
combustion products on the window. (This inlet is not in the section plane; the
intersection of the port with the main passage is circled.)
The final section of the high-pressure stack is the dilution cooler. This section houses
another perforated liner and fluid inlets. Here, high-pressure, ambient temperature water
is mixed with the combustion products to cool them below the stack BPR maximum
temperature.
The high-pressure stack is shown installed in the lab in Fig. 24. The vessel is
insulated from the black supporting stand by Macor ceramic parts. The stand is bolted to
an optical table so that optical diagnostics may be added easily.
Construction Progress
At the beginning of this project, the allocated lab space did not have adequate utilities
to run the experiment. Domestic cold water supply, lab drain, and process cooling water
lines have been installed. A 200 A, 480 V, 3 phase service was installed for the electric
heaters and triplex water pump motor. All necessary utilities upgrades are now complete.
Figure 23: Section view of high pressure stack design, showing fluid
ports, burner, axial viewing window, and liners. Ports are labeled with
their present use: TC=thermocouple, P=pressure transducer, N.C.=no
connection (these ports are plugged).
All of the systems described in the preceding section have been installed. All
pressure and flow rate sensors have been calibrated. The feed systems have been run at
ambient temperature and nominal operating pressure long enough that pressure and flow
control are well understood.
Several attempts have been made at heating the system while pressurized in
preparation for flame experiments. All of these tests were stopped after a few hours of
heating when steam was observed leaking from the high-pressure stack. It was
determined that the metal c-seals between the sections of the vessel were defective. The
seals are Inconel 718 covered with 0.001 in. silver electroplate, with a thin gold layer in
between to improve adherence of the silver. The first set of seals received had pits in the
silver layer. With most seals from this batch, there was at least one and usually several
pits on the seal line that were too large to be filled by yielding silver during installation.
The result was leaks that began small, but grew noticeably over several minutes as the
escaping water and steam eroded the leak channel. The manufacturer provided a
replacement set of seals with a smoother silver layer. These seals have operated
satisfactorily in the header section below the burner and at the windows. However, they
have continued to be problematic at the two clamped joints.
In response to the continued leaks, other sealing methods were sought. C-seals
operate by using fluid pressure to press the seal against the gland. Operating on the
theory that increased pressure could result in a good seal, a ring seal was made having a
solid cross section with a raised point on each side (Fig. 25). This seal had a diameter
and height such that it would fit in place of a c-seal. The height of this seal at its inner
diameter matched the size of the c-seal gland (0.075 in.), but the raised points were taller
than the gland height such that they were crushed during installation. It was thought that
having pressure on the seal surfaces high enough to yield stainless steel or Inconel would
provide a reliable seal. In fact, one such seal in each of stainless steel and Inconel did
Figure 24: The high-pressure stack shown without and with insulation.
seal at room temperature and during heat-up of the pressure vessel. However, they both
leaked dramatically during cool-down. It is believed that this behavior is due to thermal
expansion. During heat-up, the ring in contact with preheated water is likely warmer than
the surrounding vessel. Hence, thermal expansion keeps a positive load on the sealing
surfaces. During cool-down, the opposite thermal gradient is present, contracting the ring
faster than the vessel. This action unloads the seal and creates a large leak path. Once
these seals leaked, they would not re-seal. Either the ring had yielded to a size smaller
than the gland, or the gland had been enlarged by the warmer ring during heat-up.
The solid ring seals demonstrated that a seal could be made by applying sufficient
stress to mating surfaces. They failed due to a lack of resilience in the assembly.
Although the clamps themselves behave elastically, once the mating surfaces of the two
parts of the vessel are touching, there is nothing except fluid pressure to act on the seal.
A seal placed between the large, formerly-mating surfaces, rather than in the gland, could
remain loaded by the clamps at all times since it would hold the vessel sections apart.
However, such a seal would have to be quite thin to not have an adverse affect on the
operation of the clamps. This is because the clamps are designed to interfere with the
flanges, and too much interference will cause them to not load properly, or will prevent
assembly entirely.
After an attempt with aluminum foil, gasket seals were fabricated from 0.008 in.
fully-annealed copper sheet. One such seal is shown in Fig. 26, ready for installation.
The inner diameter is slightly smaller than the outer diameter of the gland, but is larger
than the combustor and cooler liners so as not to block the fluid flow in the annulus
outside the liner. Three locating tabs that do reach the outer diameter of the liner keep
the gasket centered during assembly. The outer diameter of the gasket is large enough
that the gaskets are strong enough to be easy to handle, but small enough that a pressure
Figure 25: Design of solid cross-section metal seal rings.
exceeding the yield stress of annealed copper is generated by the clamps.. Based on
analysis of the clamped joint and impressions on pressure-sensitive film, a pressure of
about 30,000 psi is applied to the gasket. Removed gaskets show complete impressions
of machining marks from the adjacent Inconel surfaces. This design of copper gasket has
proven successful in sealing the vessel during both heat-up and cool-down.
Fuel Choice
Even with the combustor sealed, a few more questions remained to be answered
before an attempt at flame ignition. Chief among these was whether or not the n-heptane
fuel and water would actually form a supercritical solution while flowing through the
heaters. It was important to verify this mixing since stable, steady operation of the
combustor is predicated on a stable composition and state of the reactant streams. To this
end, the stack BPR was placed near the heaters and a temporary line was run from the
fuel/water exit of the heaters to the BPR, along with a temporary dilution cooling line.
With the heater on, fuel and water flow was turned on at rates similar to what would be
used in initial experiments (20-25% mass fraction fuel, 5-10 kW firing rate). The flow
coming out of the BPR had separate phases of water and n-heptane. There was an
unpleasant, ―cooked‖ smell, indicating that some of the fuel had pyrolyzed or hydrolyzed,
but based on the volume of fuel at the outlet, most of the fuel passed through the system
in its original form. Since the fuel and water flow through tubes with an inner diameter
of 0.120 in., it is likely that there was not a sufficient interface area for the water and n-
heptane to mix thoroughly during the residence time in the heaters.
Due to the importance of fuel/water miscibility and the uncertainty of extent of
mixing, the decision was made to conduct initial experiments with a fuel that is miscible
even at ambient conditions. The fuel system was drained of n-heptane, cleaned, and
refilled with methanol. The change to methanol also required replacement of o-rings in
the solenoid shut-off valve. Additionally, a small diameter nozzle was installed in the
Figure 26: Photograph of copper foil gasket before installation
fuel/water mixing tee so that the fuel would be introduced to the water as a turbulent jet
to promote mixing.
Flame Experiments
With functioning seals and the fuel switched to methanol, the apparatus was reset for
ignition experiments. This experiment took place in December 2010. The windows were
not in place at this stage as a precaution, so indication of reaction was by temperature
data only. Figure 27 shows the methanol and oxygen mass flows and the downstream
temperature trace (measured by ―Combustor TC‖ in Fig. 23) during the first simultaneous
injection of reactants. The temperature rise indicates that exothermic reaction has
occurred. Note that the temperatures shown are subcritical since the measurement
location is about 12 in. downstream of the burner nozzle; significant cooling of the flow
occurred by this point since the vessel walls had not been completely heated to a steady-
state condition. Without optical access, it cannot yet be said that a hydrothermal flame
was produced, but it is sure that fuel and oxygen have reacted spontaneously when
brought in contact in supercritical water.
To be confident that reaction occurred, the effect of another source of the temperature
rise was tested. When reactant injection begins, the total mass flow also increases. Since
the flow is cooling a significant amount before reaching the downstream thermocouple,
Figure 27: Reactant flow rate and downstream temperature histories
during first injections of fuel and oxygen. The temperature rise after
introduction of fuel indicates that exothermic reaction has occurred.
an increased flow speed could also increase the observed temperature. After the reactant
flow was stopped (at time 181 min.), water flow continued until the temperature fell to
the pre-ignition temperature (460-470 K). Next, the water flows through the fuel/water
and oxygen/water lines were increased slightly to simulate the added mass flow from fuel
and oxygen. The effect on the temperature is much smaller than when actually injecting
fuel and oxygen; the small peak centered around 200 mins. is the result. Next, fuel and
oxygen were injected a second time to see if the first observed temperature rise was
repeatable. The temperature trace shows that exothermic reaction occurred again.
The second set of ignition experiments was conducted in January 2011. The reactant
flow and downstream temperature plots are shown in Fig. 27. This time, there is not an
obvious temperature rise accompanying the flow of methanol and oxygen except for the
second and fifth injections, indicating that reaction did not occur the other three times.
The primary parameters not shown in Figs. 27 and 28 are the temperatures of the streams
at the point of injection. These temperatures were lower during the second set of
experiments by approximately 15-20 K. Although this is a small difference compared to
the absolute temperatures, it is a large difference compared to the critical temperature.
The decreased temperatures prevailing during this set of experiments are evidently just
Figure 28: Reactant flow rate and downstream temperature histories
during the second set of flame experiments. The lack of clear
temperature increases as in Fig. 21 may indicate that the combustor was
operating just below the autoignition temperature.
below the autoignition temperature for the particular concentration of fuel that was
injected. Flow rates and stream inlet temperatures were varied throughout the second set
of experiments and only in two cases were the flows heated enough to produce a reaction.
These results indicate that the apparatus can be operated both above and below the
autoignition temperatures for various mixtures. This characteristic will allow the
mapping of autoignition and flame stability limits.
Conclusions
Experiments will transition to utilizing the sapphire windows for optical access once
the combustion process is better understood. The sapphire windows require more
controlled operation of the experiment to ensure their structural integrity, since they are
very susceptible to thermal stresses. Initial experimental results of the supercritical water
combustor demonstrate the feasibility of the use of SCWO for power generation. Further
studies of burner designs and operability limits are necessary for development of the
proposed plant.
References (Supercritical Combustor Development)
1. Wellig, B., 2003. Transpiring-wall reactor for supercritical water oxidation.
Doctoral thesis no. 15,038, Swiss Federal Institute of Technology (ETH) Zurich.
Available http://www.e-collection.ethz.ch.
Contacts (Supercritical Combustor Development)
Christopher F. Edwards: [email protected]
J.R. Heberle: [email protected]
Paul D. Mobley: [email protected]
Area 4: Aquifer Interactions. (Fendorf)
This research area involved assessing the fate of potentially toxic trace elements in
CO2-equilibrated brine solutions under elevated pressure and temperature. To
accomplish this objective, geochemical models, including EQ3/6, PHREEQC, and
MIN3P (1-3), were used to simulate post coal-combustion geochemistry of brine
solutions delivered to deep aquifers.
Bulk and Trace Element Composition of Simulated Brine and Coal
Multiple factors dictate the geochemistry of trace elements within CO2-equilibrated
brines, including 1) the original composition of the brine solution, 2) the composition of
the combusted coal, 3) the ratio of combusted coal:brine solution, 4) aquifer redox status,
and 5) trace element removal during the coal gasification/combustion process. To
establish a base-case scenario of trace element speciation and fate within a deep aquifer,
factors 1-4 (above) are explored with geochemical simulations. Factor 5 was excluded
from study since the trace elements are measured in solids (ash) removed during the
combustion process (a future task).
The composition of brine solutions may vary considerably, and are dictated by
several factors including i) the composition of porewater at the time of burial, ii) the net
effect(s) of diagenesis on porewater composition as influenced by ambient solids and
other proximal porewaters, and iii) physical transport of materials to and from a the
system via advective flow and solute mixing (4). While different combinations of these
factors may result and yield brines with disparate chemistries, sedimentary brines
typically possess salinities greater than 100 g/L and chloride concentrations greater than
10 g/L (4-6). Alkalinity may also vary considerably and range from 10 to 1000 mg/L.
For modeling purposes, a circumneutral brine solution was assumed as possessing ~1000
mmol/kg chloride and 10 mmol/kg alkalinity, along with multiple trace salts (Table 5).
Despite substantial variation, generalities can be made regarding the quantity of
potentially toxic trace elements in sub-bituminous coal. For example, trace quantities of
lead and cadmium, and to a lesser extent, arsenic and mercury, are often found. The high
coal throughput for modern power plants, however, results in large quantities of these
trace elements accruing in various combustion products, including fly ash, slag, and
gaseous emissions (7, 8). In our geochemical models, a suite of trace elements are
stipulated at concentrations representative of those found within coal combusted at
conventional plants (refs 7, 8; Table 6). Furthermore, coal used in the simulations is
assumed to possess ~67 moles of carbon/kg dry weight, which will dictate the quantity of
CO2 equilibrated with brine.
Table 6. Solid phase constituents in simulated coal combustion
Constituent mmol/kg (dw) Constituent mmol/kg (dw)
Carbon 66666 Antimony 0.04
Arsenic 0.6 Selenium 0.05
Cadmium 4.7 Zinc 0.7
Chromium 0.8 Calcium 350
Copper 0.4 Magnesium 50
Mercury 0.08 Iron 1000
Nickel 0.5 Titanium 50
Lead 1.5 Potassium 216
Sodium 600
Table 5. Aqueous Constituents in
Simulated Brine
Constituent mmol/kg H2O
Sodium 1000
Chloride 1000
Lithium 3
Potassium 4
Magnesium 30
Calcium 80
Bromide 2
Barium 2.5
Alkalinity 10
pH 7.1
Simulating the fate of trace elements in CO2-equilibrated brine solutions
The geochemical modeling software package PHREEQC in conjunction with
PITZER and WATEQ databases were used to simulate trace elements post-combustion in
brine solutions (2). PHREEQC was written to simulate the equilibrium chemistry of
aqueous solutions interacting with minerals and mineral surfaces and gas
dissolution/exosolution, and using the PITZER database allows for calculating the
activity of charged aqueous species under high ionic strength conditions (e.g. > 1M)
representative of saline brines. For geochemical species absent from the PITZER
database, the WATEQ database was used with an extended Debye-Huckle formulation to
calculate activity coefficients.
To demonstrate potential trace elements concentrations within post-combustion brine
solutions, stoichiometric concentrations of coal constituents were added incrementally to
one kg of brine, thus representing various ratios of coal to brine within the combustion
process. Reaction constituents were then free to react with the aqueous milieu, and a
variety of geochemical parameters were calculated, including the formation of aqueous
species (inclusive of ion pairs), solid phases (based on saturation indices), and gases.
Intuitively, as trace element mixture is added to one kg of brine, aqueous concentrations
of all trace elements increase (Fig. 29). However, when solid phases are allowed to
precipitate, the activity of multiple trace elements is repressed. For example, when
anoxic, low redox potential (Eh) aquifer conditions are specified, sulfate reduction is
thermodynamically favored, leading to the precipitation of trace element-bearing sulfide
minerals such as realgar, covellite, cinnabar, and galena (Fig. 30).
Thus, anoxic conditions may lower trace element concentrations (Fig. 31) owing to the
sink provided by sulfidic minerals (9).
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio
Figure 29: PHREEQC simulations of aqueous trace elements after coal
combustion and equilibration with brine. ―Coal:Brine Ratio‖ represents
fractions of coal (kg) and associated trace elements used for equilibration
with 1 kg solution. Precipitation of solid phases not allowed.
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-5
0
5
10
15
20
Realgar (AsS)
CdS
Covellite (CuS)
Cinnabar (HgS)
Galena (PbS)
Eskolaite (Cr2O3)
Anoxic Mineral Saturation Indicies
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-15
-10
-5
0
5
10
15
20
25
Delafossite (CuFeO2)
Calomel (HgCl)
Pyromorphite (Pb5(Cl(PO4)3)
Chromite (Fe,Mg)Cr2O4
Oxic Mineral Saturation Indicies
Mineral Saturation Indicies
Coal:Brine Ratio Coal:Brine RatioReaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-5
0
5
10
15
20
Realgar (AsS)
CdS
Covellite (CuS)
Cinnabar (HgS)
Galena (PbS)
Eskolaite (Cr2O3)
Anoxic Mineral Saturation Indicies
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
SI
-15
-10
-5
0
5
10
15
20
25
Delafossite (CuFeO2)
Calomel (HgCl)
Pyromorphite (Pb5(Cl(PO4)3)
Chromite (Fe,Mg)Cr2O4
Oxic Mineral Saturation Indicies
Mineral Saturation Indicies
Coal:Brine Ratio Coal:Brine Ratio
Figure 30: Mineral saturation indices for trace element-bearing minerals under anoxic
(left) and oxic (right) conditions (values exceeding 0 suggest precipitation). Arrows
represent the nominal coal:brine used with the reactor design presented in this report.
Figure 31: Aqueous concentrations of trace elements as a
function of coal:brine under anoxic conditions without solid
phase precipitation (left) and with trace element-bearing solid
phase precipitation allowed (right).
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Anoxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Anoxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Minerals may also precipitate under oxic conditions and are capable of scavenging
trace elements (e.g., delafossite, Cu-bearing phase, and chromite, a Cr-bearing solid)
(Fig. 32). Additinally, brines that become oxygenated will favor the precipitation of
Fe(III) bearing minerals such as goethite (FeOOH), which act as strong adsorbents
(scavengers) of various trace elements (10). Alternatively, Fe-oxides such as goethite
may form in-situ if aerated solutions are injected into anoxic aquifers that have an
abundance of Fe(II). Either of these scenarios will result in a decrease in trace element
concentrations due to sequestration reactions on the iron oxide surfaces (Fig. 33). In
particular, arsenic is predicted to adsorb strongly to Fe(III) (hydr)oxides, resulting in a
dramatic drop in its aqueous concentration.
Figure 33: Aqueous concentrations of trace element
contaminants as a function of coal:brine under aerated
brine conditions without (left) and with adsorption to the
iron oxide mineral goethite (right).
No Sorption Reactions
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-8
1e-7
1e-6
1e-5
1e-4
1e-3
As Cr Cu Hg Pb Zn Cd
Sorption to Goethite Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-8
1e-7
1e-6
1e-5
1e-4
1e-3
Reaction Progress
Constituent Sorption to Precipitated Goethite
Figure 32: Aqueous trace element concentrations as a function
of coal:brine under oxic conditions without solid phase
precipitation (left) and with trace element-bearing solid phase
precipitation allowed (right).
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Mineral Precipitation Allowed
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
No Solid Phases
Reaction Progress
0.000 0.005 0.010 0.015 0.020 0.025 0.030
(M)
1e-3
1e-6
1e-9
1e-12
1e-15
1e-18
1e-21
1e-24
1e-27
As Cr Cu Hg Pb Zn Cd
Oxic Brine/Effluent Equilibration
Coal:Brine Ratio Coal:Brine Ratio
Summary
High concentrations of problematic trace elements present within CO2-equilibrated
brine solutions are likely partially mitigated by secondary precipitation/adsorption
reactions occurring within the aquifer. The extent of these reactions will depend on the
redox status (aeration) of the delivered brine, as well as the redox status and
mineralogical makeup of the receiving aquifer. Future work will involve bench-scale
testing of coal combustion product equilibration with artificial brines in representative
aquifer materials, such as calcite or pyrite-bearing sandstones or shales.
References (Aquifer Interactions)
1. Mayer, K. U.; Frind, E. O.; Blowes, D. W., Multicomponent reactive transport
modeling in variably saturated porous media using a generalized formulation for
kinetically controlled reactions. Water Resour. Res 2002, 38, (9), 1174.
2. Parkhurst, D. L.; Appelo, C. A. J., User's guide to PHREEQC (version 2)-a
computer program for speciation, reaction-path, 1D-transport, and inverse
geochemial calculations. U.S. Geol. Surv. Water Resour. Inv. Rep. 1999, 99, 99-
4259.
3. Wolery, T. J. EQ3/6, A Software Package for Geochemical Modeling of Aqueous
Systems: Package Overview and Installation Guide (Version 7.0); UCRL-MA-
110662-PT-I; Lawrence Livermore National Laboratory, Livermore, California:
1992.
4. Hanor, J. S., Physical and Chemical Controls on the Composition of Waters in
Sedimentary Basins. Marine and Petroleum Geology 1994, 11, (1), 31-45.
5. Martel, A. T.; Gibling, M. R.; Nguyen, M., Brines in the Carboniferous Sydney
Coalfield, Atlantic Canada. Applied Geochemistry 2001, 16, (1), 35-55.
6. Fontes, J. C.; Matray, J. M., Geochemistry and Origin of Formation Brines from
the Paris Basin, France .1. Brines Associated with Triassic Salts. Chemical
Geology 1993, 109, (1-4), 149-175.
7. Klein, D. H.; Andren, A. W.; Carter, J. A.; Emery, J. F.; Feldman, C.; Fulkerson,
W.; Lyon, W. S.; Ogle, J. C.; Talmi, Y.; Vanhook, R. I.; Bolton, N., Pathways of
37 Trace-Elements through Coal-Fired Power-Plant. Environmental Science &
Technology 1975, 9, (10), 973-979.
8. Querol, X.; Fernandezturiel, J. L.; Lopezsoler, A., Trace-Elements in Coal and
Their Behavior During Combustion in a Large Power-Station. Fuel 1995, 74, (3),
331-343.
9. Huertadiaz, M. A.; Morse, J. W., Pyritization of Trace-Metals in Anoxic Marine-
Sediments. Geochimica Et Cosmochimica Acta 1992, 56, (7), 2681-2702.
10. Brown, G. E.; Parks, G. A., Sorption of trace elements on mineral surfaces:
Modern perspectives from spectroscopic studies, and comments on sorption in the
marine environment. International Geology Review 2001, 43, (11), 963-1073.
Contacts (Aquifer Interactions)
Scott Fendorf: [email protected]
Ben Kocar: [email protected]
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